Aci 421-1 r-08

ACI 421.1R-08
Reported by Joint ACI-ASCE Committee 421
Guide to Shear Reinforcement
for Slabs
Copyright American Concrete Institute
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Guide to Shear Reinforcement for Slabs
First Printing
June 2008
ISBN 978-0-87031-280-9
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ACI 421.1R-08 supersedes ACI 421.1R-99 and was adopted and published June 2008.
Copyright © 2008, American Concrete Institute.
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the Architect/Engineer.
Guide to Shear Reinforcement for Slabs
Reported by Joint ACI-ASCE Committee 421
ACI 421.1R-08
Tests have established that punching shear in slabs can be effectively
resisted by reinforcement consisting of vertical rods mechanically
anchored at the top and bottom of slabs. ACI 318 sets out the principles of
design for slab shear reinforcement and makes specific reference to stirrups,
headed studs, and shearheads. This guide reviews other available types
and makes recommendations for their design. The application of these
recommendations is illustrated through numerical examples.
Keywords: column-slab connection; concrete flat plate; headed shear
studs; moment transfer; prestressed concrete; punching shear; shear
stresses; shearheads; slabs; two-way slabs.
CONTENTS
Chapter 1—Introduction and scope, p. 421.1R-2
1.1—Introduction
1.2—Scope
1.3—Evolution of practice
Chapter 2—Notation and definitions, p. 421.1R-2
2.1—Notation
2.2—Definitions
Chapter 3—Role of shear reinforcement, p. 421.1R-3
Chapter 4—Punching shear design equations,
p. 421.1R-4
4.1—Strength requirement
4.2—Calculation of factored shear stress vu
4.3—Calculation of shear strength vn
4.4—Design procedure
Chapter 5—Prestressed slabs, p. 421.1R-9
5.1—Nominal shear strength
Chapter 6—Tolerances, p. 421.1R-10
Chapter 7—Requirements for seismic-resistant
slab-column connections, p. 421.1R-10
Chapter 8—References, p. 421.1R-10
8.1—Referenced standards and reports
8.2—Cited references
Appendix A—Details of shear studs, p. 421.1R-12
A.1—Geometry of stud shear reinforcement
A.2—Stud arrangements
A.3—Stud length
Appendix B—Properties of critical sections of
general shape, p. 421.1R-13
Appendix C—Values of vc within shear-reinforced
zone, p. 421.1R-14
Simon Brown*
Amin Ghali*
James S. Lai*
Edward G. Nawy
Pinaki R. Chakrabarti Hershell Gill Mark D. Marvin Eugenio M. Santiago
William L. Gamble Neil L. Hammill*
Sami H. Megally Stanley C. Woodson
Ramez B. Gayed*
Mahmoud E. Kamara*
*
Subcommittee members who prepared this report.
The committee would like to thank David P. Gustafson for his contribution to this report.
Theodor Krauthammer*
Chair
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421.1R-2 ACI COMMITTEE REPORT
Appendix D—Design examples, p. 421.1R-17
D.1—Interior column-slab connection
D.2—Edge column-slab connection
D.3—Corner column-slab connection
D.4—Prestressed slab-column connection
CHAPTER 1—INTRODUCTION AND SCOPE
1.1—Introduction
In flat-plate floors, slab-column connections are subjected
to high shear stresses produced by the transfer of the internal
forces between the columns and the slabs. Section 11.11.3 of
ACI 318-08 allows the use of shear reinforcement for slabs
and footings in the form of bars, as in the vertical legs of
stirrups. ACI 318 emphasizes the importance of anchorage
details and accurate placement of the shear reinforcement,
especially in thin slabs. Section 11.11.5 of ACI 318-08
permits headed shear stud reinforcement conforming to
ASTM A1044/A1044M. A general procedure for evaluation
of the punching shear strength of slab-column connections is
given in Section 11.11 of ACI 318-08.
Shear reinforcement consisting of vertical rods (studs) or
the equivalent, mechanically anchored at each end, can be
used. In this report, all types of mechanically anchored shear
reinforcement are referred to as “shear stud” or “stud.” To be
fully effective, the anchorage should be capable of developing
the specified yield strength of the studs. The mechanical
anchorage can be obtained by heads or strips connected to
the studs by welding. The heads can also be formed by
forging the stud ends.
1.2—Scope
Recommendations in this guide are for the design of shear
reinforcement in slabs. The design is in accordance with
ACI 318. Numerical design examples are included.
1.3—Evolution of practice
Extensive tests (Dilger and Ghali 1981; Andrä 1981; Van
der Voet et al. 1982; Mokhtar et al. 1985; Elgabry and Ghali
1987; Mortin and Ghali 1991; Dilger and Shatila 1989; Cao
1993; Brown and Dilger 1994; Megally 1998; Birkle 2004;
Ritchie and Ghali 2005; Gayed and Ghali 2006) have
confirmed the effectiveness of mechanically anchored shear
reinforcement, such as shown in Fig. 1.1, in increasing the
strength and ductility of slab-column connections subjected
to concentric punching or punching combined with moment.
Stud assemblies consisting of either a single-head stud
attached to a steel base rail by welding (Fig. 1.1(a)) or
double-headed studs mechanically crimped into a nonstructural
steel channel (Fig. 1.1(b)) are specified in ASTM A1044/
A1044M. Figure 1.2 is a top view of a slab that shows a
typical arrangement of shear reinforcement (stirrup legs or
studs) in the vicinity of an interior column. ACI 318 requires
that the spacing g between adjacent stirrup legs or studs,
measured on the first peripheral line of shear reinforcement,
be equal to or less than 2d. Requirement for distances so and
s are given in Chapter 4.
CHAPTER 2—NOTATION AND DEFINITIONS
2.1—Notation
Ac = area of concrete of assumed critical section
Av = cross-sectional area of shear reinforcement
on one peripheral line parallel to perimeter of
column section
bo = length of perimeter of critical section
cb,ct = clear concrete cover of reinforcement to
bottom and top slab surfaces, respectively
cx,cy = size of rectangular column measured in two
orthogonal span directions
D = diameter of stud or stirrup
d = effective depth of slab; average of distances
from extreme compression fiber to centroids
of tension reinforcements running in two
orthogonal directions
db = nominal diameter of flexural reinforcing bars
fc′ = specified compressive strength of concrete
fct = average splitting tensile strength of light-
weight-aggregate concrete
fpc = average value of compressive stress in
concrete in two directions (after allowance for
all prestress losses) at centroid of cross section
Fig. 1.1—Stud assemblies conforming to ASTM A1044/
A1044M: (a) single-headed studs welded to a base rail; and
(b) double-headed studs crimped into a steel channel.
Fig. 1.2—Top view of flat plate showing arrangement of
shear reinforcement in vicinity of interior column.
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SHEAR REINFORCEMENT FOR SLABS 421.1R-3
fyt = specified yield strength of shear reinforce-
ment
g = distance between adjacent stirrup legs or studs,
measured in a parallel direction to a column
face
h = overall thickness of slab
Jc = property of assumed critical section (Eq. (4-4)),
defined by ACI 318 as “analogous to polar
moment of inertia”
Jx,Jy = property of assumed critical section of any
shape, equal to d multiplied by second
moment of perimeter about x- or y-axis,
respectively (Appendix B)
Jxy = d times product of inertia of assumed shear-
critical section about nonprincipal axes x and y
(Eq. (B-11))
l = length of segment of assumed critical section
ls = overall specified height of headed stud
assembly including anchors (Fig. 1.1, Eq. (6-1))
lx,ly = projections of assumed critical section on
principal axes x and y
lx1,ly1 = lengths of sides in x and y directions of critical
section at d/2 from column face
lx2,ly2 = lengths of sides in x and y directions of critical
section at d/2 outside outermost legs of shear
reinforcement
Mux ,Muy = factored unbalanced moments transferred
between slab and column about centroidal
principal axes x and y of assumed critical section
Mux ,Muy = factored unbalanced moment about the
centroidal nonprincipal x or y axis
MuOx ,MuOy= factored unbalanced moment about x or y axis
through column’s centroid O
n = number of studs or stirrup legs per line
running in x or y direction
s = spacing between peripheral lines of shear
reinforcement
so = spacing between first peripheral line of shear
reinforcement and column face
Vp = vertical component of all effective prestress
forces crossing the critical section
Vu = factored shear force
vc = nominal shear strength provided by concrete
in presence of shear reinforcement, psi (MPa)
vn = nominal shear strength at critical section, psi
(MPa)
vs = nominal shear strength provided by shear
reinforcement, psi (MPa)
vu = maximum shear stress due to factored forces,
psi (MPa)
x,y = coordinates of point on perimeter of shear-
critical section with respect to centroidal axes
x and y
x,y = coordinates of point on perimeter of shear-
critical section with respect to centroidal
nonprincipal axes x and y
α = distance between column face and critical
section divided by d
αs = dimensionless coefficient equal to 40, 30, and
20, for interior, edge, and corner columns,
respectively
β = ratio of long side to short side of column
cross section
βp = constant used to compute vc in prestressed slabs
γvx,γvy = factor used to determine unbalanced moment
about the axes x and y between slab and
column that is transferred by shear stress at
assumed critical section
λ = modification factor reflecting the reduced
mechanical properties of lightweight concrete,
all relative to normalweight concrete of the
same compressive strength
φ = strength reduction factor = 0.75
2.2—Definitions
drop panel—thickened structural portion of a flat slab in
the area surrounding a column, as defined in Chapter 13 of
ACI 318-08. The plan dimensions of drop panels are greater
than shear capitals. For flexural strength, ACI 318 requires
that drop panels extend in each direction from the centerline
of support a distance not less than 1/6 the span length
measured from center-to-center of supports in that direction.
ACI 318 also requires that the projection of the drop panel
below the slab be at least 1/4 the slab thickness.
flat plate—flat slab without column capitals or drop panels.
shear capital—thickened portion of the slab around the
column with plan dimensions not conforming with the ACI
318 requirements for drop panels.
shear-critical section—cross section, having depth d and
perpendicular to the plane of the slab, where shear stresses
should be evaluated. Two shear-critical sections should be
considered: 1) at d/2 from column periphery; and 2) at d/2
from the outermost peripheral line of shear reinforcement (if
provided).
stud shear reinforcement (SSR)—reinforcement
conforming to ASTM A1044/A1044M and composed of
vertical rods anchored mechanically near the bottom and top
surfaces of the slab.
unbalanced moment—sum of moments at the ends of the
columns above and below a slab-column joint.
CHAPTER 3—ROLE OF SHEAR REINFORCEMENT
Shear reinforcement is required to intercept shear cracks
and prevent them from widening. The intersection of shear
reinforcement and cracks can be anywhere over the height of
the shear reinforcement. The strain in the shear reinforcement
is highest at that intersection.
Effective anchorage is essential, and its location should be
as close as possible to the structural member’s outer surfaces.
This means that the vertical part of the shear reinforcement
should be as tall as possible to avoid the possibility of cracks
passing above or below it. When the shear reinforcement is
not as tall as possible, it may not intercept all inclined shear
cracks. Anchorage of shear reinforcement in slabs is
achieved by mechanical ends (heads), bends, and hooks.
Tests (Marti 1990) have shown, however, that movement
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occurs at the bends of shear reinforcement, at Point A of
Fig. 3.1, before the yield strength can be reached in the shear
reinforcement, causing a loss of tension. Furthermore, the
concrete within the bend in the stirrups is subjected to
stresses that could potentially exceed 0.4 times the stirrup’s
yield strength fyt, causing concrete crushing. If fyt is 60 ksi
(414 MPa), the average compressive stress on the concrete
under the bend has to reach 0.4fyt for equilibrium. Because
this high stress can crush the concrete, however, slip occurs
before the development of the full fyt in the leg of the stirrup
at its connection with the bend. These difficulties, including
the consequences of improper stirrup details, were also
discussed by others (Marti 1990; Joint ACI-ASCE
Committee 426 1974; Hawkins 1974; Hawkins et al. 1975).
The movement at the end of the vertical leg of a stirrup can
be reduced by attachment to a flexural reinforcement bar, as
shown at Point B of Fig. 3.1. The flexural reinforcing bar,
however, cannot be placed any closer to the vertical leg of
the stirrup without reducing the effective slab depth d. Flexural
reinforcing bars can provide such improvement to shear
reinforcement anchorage only if attachment and direct
contact exists at the intersection of the bars (Point B of Fig. 3.1).
Under normal construction, however, it is very difficult to
ensure such conditions for all stirrups. Thus, such support is
normally not fully effective, and the end of the vertical leg of
the stirrup can move. The amount of movement is the same
for a short or long shear-reinforcing bar. Therefore, the loss
in tension is important, and the stress is unlikely to reach
yield in short shear reinforcement (in thin slabs). These prob-
lems are largely avoided if shear reinforcement is provided
with mechanical anchorage.
CHAPTER 4—PUNCHING SHEAR
DESIGN EQUATIONS
4.1—Strength requirement
This chapter presents the design procedure of ACI 318
when stirrups or headed studs are required in the slab in the
vicinity of a column transferring moment and shear. The
equations of Sections 4.3.2 and 4.3.3 apply when stirrups
and headed studs are used, respectively.
Design of critical slab sections perpendicular to the plane
of a slab should be based on
vu ≤ φvn (4-1)
in which vu is the shear stress in the critical section caused by
the transfer, between the slab and the column, of factored
shearing force or factored shearing force combined with
moment; vn is the nominal shear strength (psi or MPa); and
φ is the strength reduction factor.
Equation (4-1) should be satisfied at a critical section
perpendicular to the plane of the slab at a distance d/2 from
the column perimeter and located so that its perimeter bo is
minimum (Fig. 4.1(a)). It should also be satisfied at a critical
section at d/2 from the outermost peripheral line of the shear
reinforcement (Fig. 4.1(b)), where d is the average of
distances from extreme compression fiber to the centroids of
the tension reinforcements running in two orthogonal
directions. Figure 4.1(a) indicates the positive directions of
the internal force Vu and moments Mux and Muy that the
column exerts on the slab.
4.2—Calculation of factored shear stress vu
ACI 318 requires that the shear stress resulting from
moment transfer by eccentricity of shear be assumed to vary
linearly about the centroid of the shear-critical section. The
shear stress distribution, expressed by Eq. (4-2), satisfies this
requirement. The maximum factored shear stress vu at a critical
section produced by the combination of factored shear force
Vu and unbalanced moments Mux and Muy is
(4-2)
The coefficients γvx and γvy are given by
(4-3)
where lx1 and ly1 are lengths of the sides in the x and y directions
of a rectangular critical section at d/2 from the column face
(Fig. 4.1(a)). Appendix B gives equations for Jx, Jy, γvx, and
γvy for a shear-critical section of any shape. For a shear-critical
section in the shape of a closed rectangle, the shear stress due
to Vu combined with Muy, ACI 318 gives Eq. (4-2) with Mux =
0 and Jy replaced by Jc, which is defined as property of assumed
critical section “analogous to polar moment of inertia.” For the
closed rectangle in Fig. 4.1(a), ACI 318 gives
(4-4)
The first term on the right-hand side of this equation is equal
to Jy; the ratio of the second term to the first is commonly less
vu
Vu
Ac
-----
γvxMuxy
Jx
-------------------
γvyMuy x
Jy
-------------------+ +=
γvx 1
1
1
2
3
--- ly1 lx1⁄+
---------------------------------–=
γvy 1
1
1
2
3
--- lx1 ly1⁄+
---------------------------------–=
⎭
⎪
⎪
⎪
⎬
⎪
⎪
⎪
⎫
Jc d
lx1
3
6
--------
ly1lx1
2
2
---------------+
lx1d
3
6
-----------+=
Fig. 3.1—Geometrical and stress conditions at bend of
shear reinforcing bar.
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than 3%. The value of vu obtained by the use of Jy in Eq. (4-2)
differs on the safe side from the value obtained with Jc.
When the centroid of the shear-critical section does not
coincide with O, the centroid of the column (Fig. 4.2(b)
and (c)), the unbalanced moment Mux or Muy about the x- or
y-axis through the centroid of shear-critical section is related
to the unbalanced moment MuOx or MuOy about the x- or y-axis
through O by
Mux = MuOx + VuyO; Muy = MuOy + Vu xO (4-5)
where (xO, yO) are the coordinates of O with respect to the
centroid of the shear-critical section along the centroidal
principal x and y axes.
For the shear-critical section in Fig. 4.2(c), the moments
about the centroidal nonprincipal axes x and y (Mux and Muy)
are equivalent to the moments about the x and y axes (Mux
and Muy) that are given by Eq. (4-6).
Mux = Muxcosθ – Muysinθ; Muy = Muxsinθ + Muycosθ (4-6)
where θ is the angle of rotation of the axes x and y to coincide
with the principal axes.
4.3—Calculation of shear strength vn
Whenever the specified compressive strength of concrete
fc′ is used in Eq. (4-7a), (4-8a), (4-9a), (4-10a), and (4-12a),
its value is in pounds per square inch; when fc′ is in MPa,
Eq. (4-7b), (4-8b), (4-9b), (4-10b) and (4-12b) are used. For
prestressed slabs, refer to Chapter 5.
4.3.1 Shear strength without shear reinforcement—For
nonprestressed slabs, the shear strength of concrete at a critical
section at d/2 from column face, where shear reinforcement
is not provided, should be the smallest of
(in.-lb units) (4-7a)
(SI units) (4-7b)
where β is the ratio of long side to short side of the column
cross section
(SI units) (4-8b)
where αs is 40 for interior columns, 30 for edge columns or
20 for corner columns, and
vn = 4λ (in.-lb units) (4-9a)
vn = λ /3 (SI units) (4-9b)
At a critical section outside the shear-reinforced zone
vn = 2λ (in.-lb units) (4-10a)
vn = λ /6 (SI units) (4-10b)
Equation (4-1) should be checked first at a critical section
at d/2 from the column face (Fig. 4.1(a)). If Eq. (4-1) is not
satisfied, shear reinforcement is required.
4.3.2 Shear strength with stirrups—ACI 318 permits the
use of stirrups as shear reinforcement when d ≥ 6 in. (152 mm),
but not less than 16 times the diameter of the stirrups. When
stirrup shear reinforcement is used, ACI 318 requires that the
maximum factored shear stress at d/2 from column face
satisfy: vu ≤ 6φ (in.-lb units) (φ /2 [SI units]). The
shear strength at a critical section within the shear-reinforced
zone should be computed by
vn = vc + vs (4-11)
vn 2
4
β
---+
⎝ ⎠
⎛ ⎞ λ fc′=
vn 2
4
β
---+
⎝ ⎠
⎛ ⎞ λ
fc′
12
---------=
vn
αsd
bo
--------- 2+
⎝ ⎠
⎛ ⎞ λ fc′=
vn
αsd
bo
--------- 2+
⎝ ⎠
⎛ ⎞ λ
fc′
12
---------=
fc′
fc′
fc′
fc′
fc′ fc′
Fig. 4.1—Critical sections for shear in slab in vicinity of
interior column. Positive directions for Vu , Mux , and Muy
are indicated.
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in which
vc = 2λ (in.-lb units) (4-12a)
vc = 0.17λ (SI units) (4-12b)
and
(4-13)
where Av is the cross-sectional area of the shear reinforce-
ment legs on one peripheral line parallel to the perimeter of
the column section, and s is the spacing between peripheral
lines of shear reinforcement.
The upper limits, permitted by ACI 318, of so and the
spacing s between the peripheral lines are
so ≤ 0.5d (4-14)
s ≤ 0.5d (4-15)
where so is the distance between the first peripheral line of
shear reinforcement and the column face. The upper limit of
so is intended to eliminate the possibility of shear failure
between the column face and the innermost peripheral line of
shear reinforcement. Similarly, the upper limit of s is to avoid
failure between consecutive peripheral lines of stirrups. A line
of stirrups too close to the column can be ineffective in
intercepting shear cracks; thus, so should not be smaller
than 0.35d.
The shear reinforcement should extend away from the
column face so that the shear stress vu at a critical section at
d/2 from outermost peripheral line of shear reinforcement
(Fig. 4.1(b) and 4.2) does not exceed φvn, where vn is
calculated using Eq. (4-10a) or (4-10b).
fc′
fc′
vs
Av fyt
bos
-----------=
Fig. 4.2—Typical arrangement of shear studs and critical sections outside shear-
reinforced zone.
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4.3.3 Shear strength with studs—Section 11.4.4 of ACI
318-08 requires that: “Stirrups and other bars or wires used
as shear reinforcement shall extend to a distance d from
extreme compression fiber and shall be developed at both
ends according to 12.13.” Test results (Dilger and Ghali
1981; Andrä 1981; Van der Voet et al. 1982; Mokhtar et al.
1985; Elgabry and Ghali 1987; Mortin and Ghali 1991;
Dilger and Shatila 1989; Cao 1993; Brown and Dilger 1994;
Megally 1998; Birkle 2004; Ritchie and Ghali 2005; Gayed
and Ghali 2006) showed that studs, with anchor heads of area
equal to 10 times the cross-sectional area of the shank,
clearly satisfied this requirement. Further, the use of the
shear device, such as that shown in Fig. 1.1, demonstrated a
higher shear capacity. Other researchers (Dyken and Kepp
1988; Gayed and Ghali 2004; McLean et al. 1990; Muller et
al. 1984; Ghali et al. 1974) successfully applied other config-
urations. Based on these results, ACI 318 permits the values
given as follows when the shear reinforcement is composed
of headed studs with mechanical anchorage capable of
developing the yield strength of the rod.
The nominal shear strength provided by the concrete in the
presence of headed shear studs, using Eq. (4-11), is taken as
vc = 3λ (in.-lb units) (4-16a)
vc = λ /4 (SI units) (4-16b)
instead of 2λ (in.-lb units) (0.17λ [SI units]).
Discussion on the design value of vc is given in Appendix C.
The nominal shear strength vn (psi or MPa) resisted by
concrete and steel in Eq. (4-11) can be taken as high as 8
(in.-lb units) (2 /3 [SI units]) instead of 6 (in.-lb
units) (0.5 [SI units]). This enables the use of thinner
slabs. Experimental data showing that the higher value of vn
can be used are included in Appendix C.
ACI 318 permits upper limits for s based on the value of
vu at the critical section at d/2 from column face
s ≤ 0.75d when (in.-lb units) (0.5 [SI units]) (4-17)
s ≤ 0.5d when (in.-lb units) (0.5 [SI units]) (4-18)
When stirrups are used, ACI 318 limits s to d/2. The higher
limit for s given by Eq. (4-17) for headed shear stud spacing
is again justified by tests (Seible et al. 1980; Andrä 1981;
Van der Voet et al. 1982; Mokhtar et al. 1985; Elgabry and
Ghali 1987; Institut für Werkstoffe im Bauwesen 1996;
Regan 1996a,b; Sherif 1996).
As mentioned in Chapter 3, a vertical branch of a stirrup is
less effective than a stud in controlling shear cracks for two
reasons: 1) the shank of the headed stud is straight over its
full length, whereas the ends of the stirrup leg are curved;
and 2) the anchor heads at the top and the bottom of the stud
ensure that the specified yield strength is provided at all
sections of the shank. In a stirrup, the specified yield strength
fc′
fc′
fc′ fc′
fc′
fc′ fc′
fc′
vu
φ
---- 6 fc′≤ fc′
vu
φ
---- 6 fc′> fc′
can be developed only over the middle portion of the vertical
legs when they are sufficiently long.
Section 11.4.2 of ACI 318-08 limits the design yield
strength for stirrups as shear reinforcement to 60,000 psi
(414 MPa). Research (Otto-Graf-Institut 1996; Regan 1996a;
Institut für Werkstoffe im Bauwesen 1996) has indicated that
the performance of higher-strength studs as shear reinforcement
in slabs is satisfactory. In this experimental work, the stud
shear reinforcement in slab-column connections reached a
yield stress higher than 72,000 psi (500 MPa) without excessive
reduction of shear resistance of concrete. Thus, when studs
are used, fyt can be as high as 72,000 psi (500 MPa). In ASTM
A1044/A1044M, the minimum specified yield strength of
headed shear studs is 51,000 psi (350 MPa) based on what
was commercially available in 2005; higher yield strengths
are expected in future versions of ASTM A1044/A1044M. ACI
318 requires conformance with ASTM A1044/A1044M;
thus, it limits fyt to 51,000 psi (350 MPa).
4.3.4 Shear capitals—Figure 4.3(a) shows a shear capital
whose purpose is to increase the shear capacity without
using shear reinforcement. The plan dimensions of the shear
capital are governed by assuming that the shear strength at d/2
from the edges of the capital is governed by Eq. (4-7) to (4-9).
This type of shear capital rarely contains reinforcement other
than the vertical bars of the column because its plan dimensions
are small; with or without reinforcement, this practice is not
Fig. 4.3—Shear capital design.
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recommended. Experiments (Megally and Ghali 2002) show
that the failure of the shear capital is accompanied by a
sudden separation of wedges ABC and DEF from the shear
capital and brittle failure of the connection. The volume of
concrete within the wedges ABC and DEF is too small to
offer significant anchorage of the reinforcement that may be
provided in the shear capital to prevent the separation of the
wedges. Analyses and finite-element studies indicate that
this type of shear capital can be unsafe with a relatively low
shear force combined with high unbalanced moment
(Megally and Ghali 2002).
The plan dimensions of the shear capital should be
sufficiently large such that the maximum shear stresses at
two critical sections (Fig. 4.3(b)) satisfy Eq. (4-1). The critical
sections are at d/2 from the column face within the shear
capital, and at d/2 outside the edges of the shear capital. At
d/2 from the column, vn is calculated by Eq. (4-7) to (4-9) in
absence of shear reinforcement. At d/2 outside the edges of
the shear capital, vn is calculated by Eq. (4-10a) or (4-10b).
The extent of the shear capital should be the same as the
extent of the shear reinforcement when it is used instead of
the shear capital.
Based on experimental data, Eligehausen (1996) and
Dilger and Ghali (1981) proposed Eq. (4-19) and (4-20),
respectively, for the shear strength at critical sections at αd
from the column faces.
(SI units) (4-19a)
(SI units) (4-20b)
At α > 4, the one-way shear strength (Eq. (4-10)) is
assumed. Accordingly, as α is increased, the shear strength
(psi or MPa) drops (Fig. 4.4(a)), while the area of the shear-
critical section increases. Figure 4.4(b) shows the variation
of the shear strength, Vn = vn(α)bod for a circular column of
diameter c, transferring shearing force without unbalanced
moment. Line AB represents Vn when vn (psi) = 4
(independent of α); this greatly overestimates Vn compared
with line ACDF or EDF calculated by Eq. (4-19) or (4-20),
respectively. Line DF represents Vn with vn (psi) = 2 .
Because within Zone A to D the variation of vn is not
established, and the increase in Vn with α is not substantial,
it is herein recommended to extend the shear capital to the
zone where vn is known to be not less than the one-way
shear strength.
As a design example, consider a circular column of diameter
c, transferring a shearing force, Vu (lb) = 6 bod, where bo
= π(c + d) = the perimeter of the critical section at d/2 from
the column face in absence of the shear capital. The shear
capital that satisfies the recommended design should have an
approximate effective depth ≥1.5d, extending such that α =
1.5(c/d) + 2. It can be verified that this design will satisfy
Eq. (4-1) at the critical sections at d/2 from the column face
and at d/2 outside the edge of the shear capital.
For further justification of the recommendations in this
section, consider the slab-column connection in Fig. 4.5(a),
with a 10 in. square column supporting a 7 in. slab with d =
6 in. Based on the potential crack AB (Fig. 4.5(a)), ACI 318
permits
φVn = φbod4 ; φVn
(a)
= (348 in.2
)φ4
To increase the strength by 50%, the design in Fig. 4.5(b)
is not recommended by the present guide. If the φVn equation
is applied to the potential crack CD (Fig. 4.5(b)), the
predicted strength would be
φVn
(b)
= (576 in.2)φ4
The present guide considers the potential failure at EF,
whose slope is any angle ≤ 45 degrees. It is obvious that the
probability of failure at EF is far greater than at CD in a
design that considers the shear strength, φVn = φVn
(b)
=
(576 in.2)φ4 . This is because: 1) EF is shorter than CD;
vn α( ) 1
1 0.25α+
------------------------
⎝ ⎠
⎛ ⎞ 4 fc′ 0.5 α 4.0< <;=
vn α( ) 1
1 0.25α+
------------------------⎝ ⎠
⎛ ⎞ fc′
3
--------- 0.5 α 4.0< <;=
vn α( ) 7.5 α–
7
-----------------⎝ ⎠
⎛ ⎞ 4 fc′ 0.5 α 4.0< <;=
vn α( ) 7.5 α–
7
-----------------
⎝ ⎠
⎛ ⎞ fc′
3
--------- 0.5 α 4.0< <;=
fc′
fc′
fc′
fc′ fc′
fc′
fc′
Fig. 4.4—Variation of: (a) vn and (b) Vn, with the distance
between the shear-critical section and the column face (= αd).
--`,``,`,`,`,`,```,`,,``,`,,```-`-`,,`,,`,`,,`---

and 2) CD crosses top and bottom flexural reinforcements
whose amounts are specified by ACI 318, while EF may not
cross any reinforcement. Although that separation of the
wedge EFG (at a shearing force < φVn
(b)) may not produce
collapse, it should not be an acceptable failure. For further
justification of recommending against the design in Fig. 4.5(b),
consider the potential crack at HI that does not intercept the
shear capital. This crack can occur due to high unbalanced
moment in a direction that produces compressive stress in
the column in the vicinity of H. This guide consistently
recommends a shear-reinforced zone of the same size by the
provision of shear reinforcement or by shear capital.
4.4—Design procedure
The values of fc′, fyt, Mux, Muy, Vu, h, and d are given. The
design of shear reinforcement can be performed by the
following steps (see design examples in Appendix D):
1. At a critical section at d/2 from column face, calculate
vu and vn by Eq. (4-2) and (4-7) to (4-9). If (vu/φ) ≤ vn, no
shear reinforcement or further check is required. If (vu/φ) >
8 (in.-lb units) (2 /3 [SI units]), the slab thickness is
not sufficient; when (vu/φ) ≤ 8 (in.-lb units) (2 /3 [SI
units]), go to Step 2;
2. When (vu/φ) ≤ 6 (in.-lb units) ( /2 [SI units]),
ACI 318 permits stirrups or headed studs. When (vu/φ) >
6 (in.-lb units) ( /2 [SI units]), ACI 318 permits only
headed studs.
Calculate the contribution of concrete vc to the shear
strength (Eq. (4-12) or (4-16)) at the critical section at d/2
from column face. The difference [(vu/φ) – vc] gives the
shear stress vs to be resisted by stirrups or headed studs;
3. Select so and s within the limitations of Eq. (4-14), (4-15),
(4-17), and (4-18), and calculate the required shear reinforce-
ment area for one peripheral line Av, by solution of Eq. (4-13).
fc′ fc′
fc′ fc′
fc′ fc′
fc′ fc′
Find the minimum number of headed studs or legs of stirrups
per peripheral line;
4. Repeat Step 1 at a trial critical section at αd from
column face to find the section where (vu/φ) ≤ 2λ (in.-lb
units) (0.17λ /2 [SI units]). No other section needs to be
checked, and s is to be maintained constant. Select the
distance between the column face and the outermost peripheral
line of shear reinforcement to be ≥ [αd – (d/2)].
The position of the critical section can be determined by
selection of the number of headed studs or stirrup legs per
line, n running in x or y direction (Fig. 4.2). For example, the
distance in the x or y direction between the column face and
the critical section is equal to so + (n – 1)s + d/2. The number
n should be ≥ 2; and
5. Arrange studs to satisfy the detailing requirements
described in Appendix A.
The trial calculations involved in the aforementioned steps
are suitable for computer use (Decon 1996).
CHAPTER 5—PRESTRESSED SLABS
5.1—Nominal shear strength
When a slab is prestressed in two directions, the shear
strength of concrete at a critical section at d/2 from the
column face where shear reinforcement is not provided, is
given by (ACI 318-08):
vn = βpλ + 0.3fpc + (in.-lb units) (5-1a)
vn = βpλ + 0.3fpc + (SI units) (5-1b)
where βp is the smaller of 3.5 and [(αsd/bo) + 1.5]; fpc is the
average value of compressive stress in the two directions
(after allowance for all prestress losses) at centroid of cross
section; and Vp is the vertical component of all effective
prestress forces crossing the critical section. Equation (5-1a)
or (5-1b) is applicable only if the following are satisfied:
1. No portion of the column cross section is closer to a
discontinuous edge than four times the slab thickness h;
2. fc′ in Eq. (5-1a) (or Eq. (5-1b)) is not taken greater than
5000 psi (34.5 MPa); and
3. fpc in each direction is not less than 125 psi (0.86 MPa),
nor taken greater than 500 psi (3.45 MPa).
If any of the aforementioned conditions are not satisfied,
the slab should be treated as nonprestressed, and Eq. (4-7) to
(4-9) apply. Within the shear-reinforced zone, vn is to be
calculated by Eq. (4-11); the equations and the design
procedure in Sections 4.3.2, 4.3.3, and 4.4 apply.
In thin slabs, Vp is small with practical tendon profiles and
the slope of the tendon is hard to control. Special care should
be exercised in computing Vp in Eq. (5-1a) or (5-1b) due to
the sensitivity of its value to the as-built tendon profile.
When it is uncertain that the actual construction will match
the design assumption, a reduced or zero value for Vp should
be used in Eq. (5-1a) or (5-1b). Section D.4 is an example of
the design of the shear reinforcement in a prestressed slab.
fc′
fc′
fc′
Vp
bod
--------
fc′
12
---------
Vp
bod
--------
Fig. 4.5—Potential shear cracks. Examples of connections:
(a) without shear capital; and (b) with shear capital.
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CHAPTER 6—TOLERANCES
Shear reinforcement, in the form of stirrups or studs, can
be ineffective if the specified distances so and s are not
controlled accurately. Tolerances for these dimensions
should not exceed ±0.5 in. (±13 mm). If this requirement is
not met, a punching shear crack can traverse the slab thickness
without intersecting the shear-reinforcing elements. Tolerance
for the distance between column face and outermost peripheral
line of studs should not exceed ±1.5 in. (±38 mm).
Tests (Dilger and Ghali 1981; Andrä 1981; Van der Voet
et al. 1982; Mokhtar et al. 1985; Elgabry and Ghali 1987;
Mortin and Ghali 1991; Dilger and Shatila 1989; Cao 1993;
Brown and Dilger 1994; Megally 1998; Birkle 2004; Ritchie
and Ghali 2005; Gayed and Ghali 2006) show that headed
studs, anchored as close as possible to the top and bottom of
slabs, are effective in resisting punching shear. The designer
should specify the overall height of the stud assemblies
having the most efficiency
ls = h – ct – cb (6-1)
where h is the thickness of the member, and ct and cb are the
specified concrete covers at top and bottom, respectively.
ACI 318 permits a manufacturing tolerance: the actual
overall height can be shorter than ls by no more than db/2,
where db is the diameter of the tensile flexural reinforcement
(Fig. 6.1). In slabs in the vicinity of columns, the tensile flexural
reinforcement is commonly at the top; in footings, the tensile
flexural reinforcement is commonly at the bottom.
CHAPTER 7—REQUIREMENTS FOR SEISMIC-
RESISTANT SLAB-COLUMN CONNECTIONS
Connections of columns with flat plates should not be
considered in design as part of the system resisting lateral
forces. Due to the lateral movement of the structure in an
earthquake, however, the slab-column connections transfer
vertical shearing force V combined with reversals of moment
M. Experiments (Cao 1993; Brown and Dilger 1994;
Megally 1998; Ritchie and Ghali 2005; Gayed and Ghali
2006) were conducted on slab-column connections to simulate
the effect of interstory drift in a flat plate structure. In these
tests, the column transferred a constant shearing force V and
cyclic moment reversals with increasing magnitude. The
experiments showed that, when the slab was provided with
shear headed stud reinforcement, the connections behaved in
a ductile fashion. They could withstand, without failure, drift
ratios that varied between 3 and 7%, depending upon the
magnitude of V. The drift ratio is defined as the difference
between the lateral displacements of two successive floors
divided by the floor height. For a given value of Vu, the slab
can resist a moment Mu, which can be determined by the
procedure and equations given in Chapter 4; the value of vc
(Eq. (4-12) or (4-16)), however, should be limited to
vc = 1.5 (in.-lb units) (7-1a)
vc = /8 (SI units) (7-1b)
fc′
fc′
This reduced value of vc is based on the experiments
mentioned in this section, which indicate that the concrete
contribution to the shear resistance is diminished by the
moment reversals. This reduction is analogous to the reduction
of vc to 0 that is required by ACI 318 for framed members.
ACI 421.2R gives recommendations for designing flat plate-
column connections with sufficient ductility to go through
lateral drift due to earthquakes without punching shear
failure or loss of moment transfer capacity. A report on tests
at the University of Washington (Hawkins 1984) does not
recommend the aforementioned reduction of vc (Eq. (7-1)).
CHAPTER 8—REFERENCES
8.1—Referenced standards and reports
The documents of the various standards-producing organi-
zations, referred to in this document, are listed below with
their serial designations.
318 Building Code Requirements for Structural Concrete
421.2R Seismic Design of Punching Shear Reinforcement
in Flat Plates
ASTM International
A1044/ Specification for Steel Stud Assemblies for Shear
A1044M Reinforcement of Concrete
Canadian Standards Association
A23.3 Design of Concrete Structures for Buildings
The above publications may be obtained from the
following organizations:
P.O. Box 9094
Farmington Hills, MI 48333-9094
www.concrete.org
ASTM International
100 Barr Harbor Dr.
West Conshohocken, PA 19428-2959
www.astm.org
Fig. 6.1—Section in slab perpendicular to shear stud line.
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Canadian Standards Association
178 Rexdale Blvd.
Rexdale, Ontario M9W 1R3
Canada
www.csa.ca
8.2—Cited references
Andrä, H. P., 1981, “Strength of Flat Slabs Reinforced with
Stud Rails in the Vicinity of the Supports (Zum Tragverhalten
von Flachdecken mit Dübelleisten—Bewehrung im Auflager-
bereich),” Beton und Stahlbetonbau, Berlin, V. 76, No. 3, Mar.,
pp. 53-57, and No. 4, Apr., pp. 100-104.
Birkle, G., 2004, “Punching of Slabs: Thickness and Stud
Layout,” PhD dissertation, University of Calgary, Calgary,
AB, Canada, 152 pp.
Brown, S., and Dilger, W. H., 1994, “Seismic Response of
Flat-Plate Column Connections,” Proceedings, Canadian
Society for Civil Engineering Conference, V. 2, Winnipeg,
MB, Canada, pp. 388-397.
Cao, H., 1993, “Seismic Design of Slab-Column
Connections,” MSc thesis, University of Calgary, Calgary,
AB, Canada, 188 pp,
Decon, 1996, “STDESIGN,” Computer Program for
Design of Shear Reinforcement for Slabs, Decon, Brampton,
ON, Canada.
Dilger, W. H., and Ghali, A., 1981, “Shear Reinforcement
for Concrete Slabs,” Proceedings, ASCE, V. 107, No. ST12,
Dec., pp. 2403-2420.
Dilger, W. H., and Shatila, M., 1989, “Shear Strength of
Prestressed Concrete Edge Slab-Columns Connections with
and without Stud Shear Reinforcement,” Canadian Journal
of Civil Engineering, V. 16, No. 6, pp. 807-819.
Dyken, T., and Kepp, B., 1988, “Properties of T-Headed
Reinforcing Bars in High-Strength Concrete,” Publication
No. 7, Nordic Concrete Research, Norske Betongforening,
Oslo, Norway, Dec.
Elgabry, A. A., and Ghali, A., 1987, “Tests on Concrete
Slab-Column Connections with Stud Shear Reinforcement
Subjected to Shear-Moment Transfer,” ACI Structural
Journal, V. 84, No. 5, Sept.-Oct., pp. 433-442.
Elgabry, A. A., and Ghali, A., 1996, “Moment Transfer by
Shear in Slab-Column Connections,” ACI Structural
Journal, V. 93, No. 2, Mar.-Apr., pp. 187-196.
Eligehausen, R., 1996, “Bericht über Zugversuche mit
Deha Kopfbolzen (Report on Pull Tests on Deha Anchor
Bolts),” Institut für Werkstoffe im Bauwesen, University of
Stuttgart, Report No. DE003/01-96/32, Sept. (Research
carried out on behalf of Deha Ankersystene, GMBH & Co.,
Gross-Gerau, Germany)
Gayed, R. B., and Ghali, A., 2004, “Double-Head Studs as
Shear Reinforcement in Concrete I-Beams,” ACI Structural
Journal, V. 101, No. 4, July-Aug., pp. 549-557.
Gayed, R. B., and Ghali, A., 2006, “Seismic-Resistant
Joints of Interior Columns with Prestressed Slabs,” ACI
Structural Journal, V. 103, No. 5, Sept.-Oct., pp. 710-719.
Also see Errata in ACI Structural Journal, V. 103, No. 6,
Nov.-Dec. 2006, p. 909.
Ghali, A.; Sargious, M. A.; and Huizer, A., 1974, “Vertical
Prestressing of Flat Plates around Columns,” Shear in
Reinforced Concrete, SP-42, American Concrete Institute,
Farmington Hills, MI, pp. 905-920.
Hawkins, N. M., 1974, “Shear Strength of Slabs with Shear
Reinforcement,” Shear in Reinforced Concrete, SP-42, Amer-
ican Concrete Institute, Farmington Hills, MI, pp. 785-815.
Hawkins, N. M., 1984, “Response of Flat Plate Concrete
Structures to Seismic and Wind Forces,” Report SM84-1,
University of Washington, July.
Hawkins, N. M.; Mitchell, D.; and Hanna, S. H., 1975,
“The Effects of Shear Reinforcement on Reversed Cyclic
Loading Behavior of Flat Plate Structures,” Canadian
Journal of Civil Engineering, V. 2, No. 4, Dec., pp. 572-582.
Hoff, G. C., 1990, “High-Strength Lightweight Aggregate
Concrete—Current Status and Future Needs,” Proceedings,
2nd International Symposium on Utilization of High-
Strength Concrete, Berkeley, CA, May, pp. 20-23.
Institut für Werkstoffe im Bauwesen, 1996, “Bericht über
Versuche an punktgestützten Platten bewehrt mit DEHA
Doppelkopfbolzen und mit Dübelleisten (Test Report on
Point Supported Slabs Reinforced with DEHA Double Head
Studs and Studrails),” Universität—Stuttgart, Report No. AF
96/6 – 402/1, Germany, DEHA, 81 pp.
Joint ACI-ASCE Committee 426, 1974, “The Shear
Strength of Reinforced Concrete Members—Slabs,” Journal
of the Structural Division, ASCE, V. 100, No. ST8, Aug.,
pp. 1543-1591.
Leonhardt, F., and Walther, R., 1965, “Welded Wire Mesh
as Stirrup Reinforcement: Shear on T-Beams and Anchorage
Tests,” Bautechnik, V. 42, Oct. (in German)
Mart, P.; Parlong, J.; and Thurlimann, B., 1977, “Schub-
versuche and Stahlbeton-Platten,” Bericht Nr. 7305-2,
Institut fur Baustatik aund Konstruktion, ETH Zurich,
Birkhauser Verlag, Basel and Stuttgart, Germany.
Marti, P., 1990, “Design of Concrete Slabs for Transverse
Shear,” ACI Structural Journal, V. 87, No. 2, Mar.-Apr.,
pp. 180-190.
McLean, D.; Phan, L. T.; Lew, H. S.; and White, R. N.,
1990, “Punching Shear Behavior of Lightweight Concrete
Slabs and Shells,” ACI Structural Journal, V. 87, No. 4,
July-Aug., pp. 386-392.
Megally, S. H., 1998, “Punching Shear Resistance of
Concrete Slabs to Gravity and Earthquake Forces,” PhD
dissertation, University of Calgary, Calgary, AB, Canada,
468 pp.
Megally, S. H., and Ghali, A., 1996, “Nonlinear Analysis of
Moment Transfer between Columns and Slabs,” Proceedings,
V. IIa, Canadian Society for Civil Engineering Conference,
Edmonton, AB, Canada, pp. 321-332.
Megally, S. H., and Ghali, A., 2002, “Cautionary Note on
Shear Capitals,” Concrete International, V. 24, No. 3, Mar.,
pp. 75-83.
Mokhtar, A. S.; Ghali, A.; and Dilger, W. H., 1985, “Stud
Shear Reinforcement for Flat Concrete Plates,” ACI JOURNAL,
Proceedings V. 82, No. 5, Sept.-Oct., pp. 676-683.
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Mortin, J., and Ghali, A., 1991, “Connection of Flat Plates
to Edge Columns,” ACI Structural Journal, V. 88, No. 2,
Mar.-Apr., pp. 191-198.
Muller, F. X.; Muttoni, A.; and Thurlimann, B., 1984,
“Durchstanz Versuche an Flachdecken mit Aussparungen
(Punching Tests on Slabs with Openings),” ETH Zurich,
Research Report No. 7305-5, Birkhauser Verlag, Basel and
Stuttgart, Germany.
Otto-Graf-Institut, 1996, “Durchstanzversuche an Stahl-
betonplatten mit Rippendübeln und Vorgefertigten Gross-
flächentafeln (Punching Shear Tests on Concrete Slabs with
Deformed Studs and Large Precast Slabs),” Report No. 21-
21634, University of Stuttgart, Germany, July.
Regan, P. E., 1996a, “Double Headed Studs as Shear
Reinforcement—Tests of Slabs and Anchorages,” University
of Westminster, London, England, Aug.
Regan, P. E., 1996b, “Punching Test of Slabs with Shear
Reinforcement,” University of Westminster, London,
England, Nov.
Ritchie, M., and Ghali, A., 2005, “Seismic-Resistant
Connections of Edge Columns with Prestressed Slabs,” ACI
Structural Journal, V. 102, No. 2, Mar.-Apr., pp. 314-323.
Seible, F.; Ghali, A.; and Dilger, W. H., 1980, “Preassembled
Shear Reinforcing Units for Flat Plates,” ACI JOURNAL,
Proceedings V. 77, No. 1, Jan.-Feb., pp. 28-35.
Sherif, A., 1996, “Behavior of R.C. Flat Slabs,” PhD
dissertation, University of Calgary, Calgary, AB, Canada,
397 pp.
Van der Voet, F.; Dilger, W. H.; and Ghali, A., 1982,
“ConcreteFlatPlateswithWell-Anchored Shear Reinforcement
Elements,” Canadian Journal of Civil Engineering, V. 9,
No. 1, pp. 107-114.
APPENDIX A—DETAILS OF SHEAR STUDS
A.1—Geometry of stud shear reinforcement
Several types and configurations of shear studs have been
reported in the literature. Shear studs mounted on a continuous
steel strip, as discussed in the main text of this report, have
been developed and investigated (Dilger and Ghali 1981;
Andrä 1981; Van der Voet et al. 1982; Mokhtar et al. 1985;
Elgabry and Ghali 1987; Mortin and Ghali 1991; Dilger and
Shatila 1989; Cao 1993; Brown and Dilger 1994; Megally
1998; Birkle 2004; Ritchie and Ghali 2005; Gayed and Ghali
2006). Headed reinforcing bars were developed and applied
in Norway (Dyken and Kepp 1988) for high-strength
concrete structures, and it was reported that such applications
improved the structural performance significantly (Gayed
and Ghali 2004; Hoff 1990). Another type of headed shear
reinforcement was implemented for increasing the punching
shear strength of lightweight concrete slabs and shells
(McLean et al. 1990). Several other approaches for mechanical
anchorage in shear reinforcement can be used (Marti 1990;
Muller et al. 1984; Mart et al. 1977; Ghali et al. 1974).
Several types are depicted in Fig. A.1. ACI 318 permits stirrups
in slabs with d ≥ 6 in. (152 mm), but not less than 16 times
the diameter of the stirrups. In the stirrup details shown in
Fig. A.1(a) (from ACI 318), a bar has to be lodged in each
bend to provide the mechanical anchorage necessary for the
development of fyt in the vertical legs. Matching this detail
and the design spacing so and s in actual construction ensure
the effectiveness of stirrups as assumed in design.
The anchors should be in the form of circular or rectan-
gular plates, and their area should be sufficient to develop the
specified yield strength of studs, fyt. ASTM A1044/A1044M
specifies an anchor head area equal to 10 times the cross-
sectional area of the stud. It is recommended that the
performance of the shear stud reinforcement be verified
before their use.
A.2—Stud arrangements
Shear studs in the vicinity of rectangular columns should
be arranged on peripheral lines. The term “peripheral line” is
used in this report to mean a line running parallel to and at
constant distance from the sides of the column cross section.
Figure 4.2 shows a typical arrangement of stud shear reinforce-
ment in the vicinity of a rectangular interior, edge, and corner
columns. Tests (Dilger and Ghali 1981) showed that studs
are most effective near column corners. For this reason,
shear studs in Fig. 4.2 are aligned with column faces. In the
direction parallel to a column face, the distance g between
Fig. A.1—Shear reinforcement Type (a) is copied from ACI
318. Types (b) to (e) are from Dyken and Kepp (1988),
Gayed and Ghali (2004), McLean et al. (1990), Muller et al.
(1984), and Ghali et al. (1974).
--`,``,`,`,`,`,```,`,,``,`,,```-`-`,,`,,`,`,,`---

lines of shear studs should not exceed 2d, where d is the
effective depth of the slab. When stirrups are used, the same
limit for g should be observed (Fig. A.1(a)).
The stud arrangement for circular columns is shown in
Fig. A.2. The minimum number of peripheral lines of shear
studs, in the vicinity of rectangular and circular columns, is two.
A.3—Stud length
The studs are most effective when their anchors are as
close as possible to the top and bottom surfaces of the slab.
Unless otherwise protected, the minimum concrete cover of
the anchors should be as required by ACI 318. The cover of
the anchors should not exceed the minimum cover plus
one-half bar diameter of flexural reinforcement (Fig. 6.1).
The mechanical anchors should be placed in the forms above
reinforcement supports, which ensure the specified concrete
cover.
APPENDIX B—PROPERTIES OF CRITICAL
SECTIONS OF GENERAL SHAPE
Figure B.1 shows the top view of critical sections for shear
in slabs. The centroidal principal x and y axes of the critical
sections, Vu, Mux, and Muy are shown in their positive
directions. The shear force Vu acts at the column centroid;
Vu, Mux, and Muy represent the effects of the column on the
slab. lx and ly are projections of the shear-critical sections on
directions of principal x and y axes.
The coefficients γvx and γvy are given by Eq. (B-1) to (B-6).
ACI 318-08 gives Eq. (B-1) and (B-2); Eq. (B-3) to (B-6) are
based on finite-element studies (Elgabry and Ghali 1996;
Megally and Ghali 1996).
Interior column-slab connections (Fig. B.1(a))
(B-1)γvx 1
1
1
2
3
--- ly lx⁄+
----------------------------–=
Fig. A.2—Shear headed stud reinforcement arrangement for
circular columns.
Fig. B.1—Shear-critical sections outside shear-reinforced
zones and sign convention of factored internal forces trans-
ferred from columns to slabs.
--`,``,`,`,`,`,```,`,,``,`,,```-`-`,,`,,`,`,,`---

(B-2)
Edge column-slab connections (Fig. B.1(b))
(B-3)
but γvy = 0 when < 0.2 (B-4)
Corner column-slab connections (Fig. B.1(c))
γvx = 0.4 (B-5)
but γvy = 0 when < 0.2 (B-6)
Equations (B-7) to (B-9) give the values of Ac, Jx, and Jy
that determine by Eq. (4-2) the distribution of shear stress vu,
whose resultant components are exactly Vu, γvx Mux, and
γvyMuy. Generally, the critical section perimeter can be
considered as composed of straight segments. The values of
Ac, Jx, and Jy can be determined by summation of the contri-
bution of the segments
(B-7)
(B-8)
(B-9)
where xi, yi, xj, and yj are coordinates of points i and j at the
extremities of a typical segment whose length is l. For a
circular shear-critical section, Ac = 2πd (radius) and Jx = Jy
= πd (radius)3
.
When the critical section has no axis of symmetry, such as
in Fig. 4.2(c), the centroidal principal axes can be deter-
mined by the rotation of the centroidal nonprincipal x and y
axes an angle θ, given by
(B-10)
The absolute value of θ is less than π/2; when the value is
positive, θ is measured in the clockwise direction. Jx and Jy
can be calculated by Eq. (B-8) and (B-9), substituting x and
y for x and y. Jxy is equal to d times the product of inertia of
the perimeter of the critical section about the centroidal
nonprincipal x and y axes
(B-11)
The coordinates of any point on the perimeter of the critical
section with respect to the centroidal principal axes can be
calculated by Eq. (B-12) and (B-13)
x = xcosθ + ysinθ (B-12)
y = –xsinθ + ycosθ (B-13)
The xand y coordinates, determined by Eq. (B-12) and (B-13),
can now be substituted in Eq. (B-8) and (B-9) to give the
values of Jx and Jy.
When the maximum vu occurs at a single point on the critical
section, rather than on a side, the peak value of vu does not
govern the strength due to stress redistribution (Brown and
Dilger 1994). In this case, vu may be investigated at a point
located at a distance 0.4d from the peak point. This will give
a reduced vu value compared with the peak value; the reduction
should not be allowed to exceed 15%.
APPENDIX C—VALUES OF vc WITHIN
SHEAR-REINFORCED ZONE
This design procedure of the shear reinforcement requires
calculation of vn = vc + vs at the critical section at d/2 from
the column face. The value allowed for vc is 2 (in.-lb
units) ( /6 [SI units]) when stirrups are used, and 3
(in.-lb units) ( /4 [SI units]) when headed shear studs are
used. The reason for the higher value of vc for slabs with
headed shear stud reinforcement is the almost slip-free
anchorage of the studs. In structural elements reinforced with
conventional stirrups, the anchorage by hooks or 90-degree
bends is subject to slip, which can be as high as 0.04 in. (1 mm)
when the stress in the stirrup leg approaches its yield strength
(Leonhardt and Walther 1965). This slip is detrimental to the
effectiveness of stirrups in slabs because of their relative
small depth compared with beams. The influence of the slip
is manifold:
• Increase in width of the shear crack;
• Extension of the shear crack into the compression zone;
• Reduction of the shear resistance of the compression
zone; and
• Reduction of the shear friction across the crack.
All of these effects reduce the shear capacity of the
concrete in slabs with stirrups. To reflect the stirrup slip in
the shear resistance equations, refinement of the shear failure
model is required. The empirical equation vn = vc + vs,
adopted in almost all codes, is not the ideal approach to solve
the shear design problem. A mechanics-based model that is
acceptable for codes is not presently available. There is,
however, enough experimental evidence that use of the
empirical equation vn = vc + vs with vc = 3 (in.-lb units)
γvy 1
1
1
2
3
--- lx ly⁄+
----------------------------–=
γvx 1
1
1
2
3
--- ly lx⁄+
----------------------------–=
γvy 1
1
1
2
3
---
lx
ly
--- 0.2–+
----------------------------------–=
lx
ly
---
γvy 1 1
1 2
3
---
lx
ly
--- 0.2–+
----------------------------------–=
lx
ly
---
Ac d l∑=
Jx d l
3
--- yi
2
yiyj yj
2
+ +( )∑=
Jx d
l
3
--- xi
2
xixj xj
2
+ +( )∑=
2θtan
2Jxy
–
Jx Jy–
---------------=
Jxy
d
l
6
--- 2xiyi xiyj xjyi 2xjyj+ + +( )∑=
fc′
fc′ fc′
fc′
fc′
--`,``,`,`,`,`,```,`,,``,`,,```-`-`,,`,,`,`,,`---

( /4 [SI units]) gives a safe design for slabs with shear
headed stud reinforcement. This approach is adopted in the
Canadian code, CSA A23.3.
Numerous test slab-column connections reinforced with
headed studs are reported in the literature (Table C.1). In the
majority of these tests, the failure is at sections outside the
shear-reinforced zone. Table C.2 lists only the tests in which
the failure occurred within the shear-reinforced zone.
Column 12 of Table C.2 gives the ratio vtest/vcode, where
vcode is the value allowed by ACI 318, with vc = 3 (in.-
lb units) ( /4 [SI units]) (Eq. (4-16a) or (4-16b)). The
values of vtest/vcode greater than 1.0 indicate there is safety
of design with vc = 3 (in.-lb units) ( /4 [SI units]).
Table C.3 summarizes experimental data of numerous
slabs in which the maximum shear stress vu obtained in test,
at the critical section at d/2 from column face, reaches or
exceeds 8 (in.-lb units) (2 /3 [SI units]). Table C.3
indicates that vn can be safely taken equal to 8 (in.-lb
units) (2 /3 [SI units]) (Section 4.3.3).
Table C.4 gives the experimental results of slabs having
stud shear reinforcement with the spacing between headed
studs greater or close to the upper limit given by Eq. (4-17).
In Table C.4, vcode is the nominal shear stress calculated by
ACI 318, with the provisions given in Section 4.3.3. The
value vcode is calculated at d/2 from column face when
failure is within the shear-reinforced zone, or at a section at
d/2 from the outermost studs when failure occurs outside the
shear-reinforced zone. The ratio vtest/vcode greater than 1.0
indicates that it is safe to use headed studs spaced at the
upper limit set by Eq. (4-17) and to calculate the strength
with the provisions in Section 4.3.3.
fc′
fc′
fc′
fc′ fc′
fc′ fc′
fc′
fc′
Table C.1—List of references on slab-column connections tests using stud shear reinforcement
Experiment no. Reference Experiment no. Reference Experiment no. Reference
1 to 5 Andrä 1981 16 to 18 Regan 1996a 26 to 29 Elgabry and Ghali 1987
6, 7 Footnote* 19, 20 Regan 1996b 30 to 36 Mokhtar et al. 1985
8, 9 Otto-Graf-Institut 1996 21 to 24, 37 Sherif 1996 42,43 Seible et al. 1980
10 to 15 Intitut für Werkstoffe im
Bauwesen 1996 25, 38 to 41 Van der Voet et al. 1982 — —
*
“Grenzzustände der Tragfäkigheit für Durchstanzen von Platten mit Dübelleistein nach EC2 (Ultimate Limit States of Punching of Slabs with Studrails According to EC2),” Stuttgart,
Germany, 1996, 15 pp.
Table C.2—Slabs with stud shear reinforcement failing within shear-reinforced zone
Experiment
Square
column size,
in. (mm)
fc′, psi
(MPa)
d, in.
(mm) s/d
Tested capacities
M at critical
section centroid,
kip-in. (kN-m)
Maximum
shear stress
vu, psi (MPa)
fyt , ksi
(MPa)
Av, in.2
(mm2) vtest /vcode
†
Remarks
Vu, kips
(kN)
Mu, kip-in.
(kN-m)
(1) (2) (3) (4) (5) (6) (7) (8) (9) (10) (11) (12) (13)
20
7.87
(200)
5660
(39.0)
6.30
(160) 0.75
214
(952) 0 0
599
(4.13)
64.1
(442)
1.402
(905) 1.00 Interior column
21 9.84
(250)
4100
(28.3)
4.49
(114)
0.70 47.4
(211)
651
(73.6)
491
(55.5)
528
(3.64)
55.1
(380)
0.66
(426)
1.14 Edge column
22 9.84
(250)
4030
(27.8)
4.49
(114) 0.70 52.8
(235)
730
(82.5)
552
(62.4)
590
(4.07)
55.1
(380)
0.66
(426) 1.28 Edge column
23 9.84
(250)
4080
(28.1)
4.49
(114) 0.70 26.0
(116)
798
(90.2)
708
(80.0)
641
(4.42)
55.1
(380)
0.66
24 9.84
(250)
4470
(30.8)
4.49
(114) 0.70 27.2
(121)
847
(95.7)
755
(85.3)
693
(4.78)
55.1
(380)
0.66
26 9.84
(250)
4890
(33.7)
4.49
(114) 0.75 34.0
(151)
1434
(162.0)
1434
(162.0)
570
(3.93)
66.7
(460)
1.570
27
9.84
(250)
5660
(39.0)
4.49
(114) 0.75
67.0
(298)
1257
(142.0)
1257
(142.0)
641
(4.42)
66.7
(460)
1.570
28 9.84
(250)
5920
(40.8)
4.49
(114)
0.5 and
0.95
67.0
(298)
1328
(150.1)
1328
(150.1)
665
(4.59)
66.7
(460)
0.880
(568)
1.08 Interior column
29 9.84
(250)
6610
(45.6)
4.49
(114)
0.5 and
0.97
101
(449)
929
(105)
929
(105)
673
(4.64)
66.7
(460)
0.880
30* 9.84
(250)
5470
(37.7)
4.49
(114) 0.75 117
(520) 0 0 454
(3.13)
40.3
(278)
1.320
39 9.84
(250)
4210
(29.0)
4.49
(114) 0.88 113
(507) 0 0 444
(3.06)
47.1
(325)
0.460
Mean 1.18
Coefficient of variation 0.17
*
Semi-lightweight concrete; is replaced in calculation by fct/6.7; fct is average splitting tensile strength of lightweight aggregate concrete; fct used herein = 377 psi (2.60 MPa),
determined experimentally.
†
vcode is smaller of 8 , psi (2 /3, MPa) and (3 + vs, psi) ( /4 + vs, MPa), where vs = Av fyt /(bos).
fc′
fc′ fc′ fc′ fc′
--`,``,`,`,`,`,```,`,,``,`,,```-`-`,,`,,`,`,,`---

Table C.3—Tests with maximum vu at critical section of d/2 from column face exceeding 8 psi (2 /3
MPa) (slabs with stud shear reinforcement)
Experiment
Column size, in.
(mm)*
fc′, psi
(MPa)
8 , psi
(2/ 3, MPa)
Tested capacities
d, in. (mm)
M at critical
section centroid,
kip-in. (kN-m)
Maximum shear
stress vu, psi
(MPa) vu/8V, kips (kN)
M, kip-in.
(kN-m)
(1) (2) (3) (4) (5) (6) (7) (8) (9) (10)
1 11.81 sq. (300 sq.) 6020 (41.5) 621 (4.28) 476 (2120) 0 9.06 (230) 0 629 (4.24) 1.07
2 11.81 sq. (300 sq.) 5550 (38.3) 589 (4.06) 428 (1900) 0 8.86 (225) 0 585 (4.03) 1.00
3 11.81 sq. (300 sq.) 3250 (22.4) 456 (3.14) 346 (1540) 0 8.66 (220) 0 488 (3.37) 1.07
4 19.68 cr. (500 cr.) 5550 (38.3) 589 (4.06) 665 (2960) 0 10.51 (267) 0 667 (4.60) 1.13
5 14.57 sq. (370 sq.) 6620 (45.7) 651 (4.49) 790 (3510) 0 11.22 (285) 0 682 (4.70) 1.05
6 12.60 cr. (320 cr.) 5870 (40.5) 613 (4.23) 600 (2670) 0 9.33 (237) 0 934 (6.44) 1.52
7 12.60 cr. (320 cr.) 6020 (41.5) 621 (4.28) 620 (2760) 0 9.33 (237) 0 965 (6.66) 1.55
8 10.23 sq. (260 sq.) 3120 (21.5) 447 (3.08) 271 (1200) 0 8.07 (205) 0 459 (3.17) 1.03
9 10.23 sq. (260 sq.) 3270 (22.6) 457 (3.15) 343 (1530) 0 8.07 (205) 0 582 (4.01) 1.27
10 7.48 cr. (190 cr.) 3310 (22.8) 460 (3.17) 142 (632) 0 5.83 (148) 0 582 (4.01) 1.26
11 7.48 cr. (190 cr.) 3260 (22.5) 456 (3.14) 350 (1560) 0 9.60 (244) 0 679 (4.68) 1.48
12 7.48 cr. (190 cr.) 4610 (31.8) 543 (3.74) 159 (707) 0 6.02 (153) 0 623 (4.30) 1.14
13 7.48 cr. (190 cr.) 3050 (21.0) 441 (3.04) 128 (569) 0 5.91 (150) 0 516 (3.56) 1.17
14 7.48 cr. (190 cr.) 3340 (23.0) 462 (3.19) 278 (1240) 0 9.72 (247) 0 530 (3.66) 1.14
15 7.48 cr. (190 cr.) 3160 (21.8) 449 (3.10) 255 (1130) 0 9.76 (248) 0 482 (3.32) 1.07
16 9.25 cr. (235 cr.) 4630 (31.9) 544 (3.75) 207 (921) 0 5.94 (151) 0 728 (5.02) 1.34
17 9.25 cr. (235 cr.) 5250 (36.2) 580 (4.00) 216 (961) 0 6.14 (156) 0 725 (5.00) 1.25
18 9.25 cr. (235 cr.) 5290 (36.5) 582 (4.01) 234 (1040) 0 6.50 (165) 0 725 (5.00) 1.24
19 7.87 sq. (200 sq.) 5060 (34.9) 569 (3.92) 236 (1050) 0 6.30 (160) 0 661 (4.56) 1.16
20 7.87 sq. (200 sq.) 5660 (39.0) 601 (4.14) 214 (952) 0 6.30 (160) 0 599 (4.13) 1.00
21† 9.84 sq. (250 sq.) 4100 (28.3) 513 (3.54) 47.4 (211) 651 (73.6) 4.49 (114) 491 (55.5) 528 (3.64) 1.03
22† 9.84 sq. (250 sq.) 4030 (27.8) 508 (3.50) 52.8 (235) 730 (82.5) 4.49 (114) 552 (62.4) 590 (4.07) 1.16
23† 9.84 sq. (250 sq.) 4080 (28.1) 511 (3.52) 26.9 (120) 798 (90.2) 4.49 (114) 708 (80.0) 641 (4.42) 1.25
24† 9.84 sq. (250 sq.) 4470 (30.8) 535 (3.69) 27.2 (121) 847 (95.7) 4.49 (114) 755 (85.3) 693 (4.78) 1.29
25 9.84 sq. (250 sq.) 4280 (29.5) 523 (3.61) 135 (600) 0 4.45 (113) 0 532 (3.67) 1.02
26 9.84 sq. (250 sq.) 4890 (33.7) 559 (3.86) 33.7 (150) 1434 (162.0) 4.49 (114) 1434 (162.0) 570 (3.93) 1.02
27 9.84 sq. (250 sq.) 5660 (39.0) 602 (4.15) 67.4 (300) 1257 (142.0) 4.49 (114) 1257 (142.0) 641 (4.42) 1.06
28 9.84 sq. (250 sq.) 5920 (40.8) 615 (4.24) 67.4 (300) 1328 (150.1) 4.49 (114) 1328 (150.1) 665 (4.59) 1.08
29 9.84 sq. (250 sq.) 6610 (45.6) 651 (4.49) 101 (449) 929 (105) 4.49 (114) 924 (104) 673 (4.64) 1.03
Mean 1.17
*
Column 2 gives side dimension of square (sq.) columns or diameter of circular (cr.) columns.
†
Edge slab-column connections. Other experiments are on interior slab-column connections.
fc′ fc′
fc′
fc′ fc′
--`,``,`,`,`,`,```,`,,``,`,,```-`-`,,`,,`,`,,`---

APPENDIX D—DESIGN EXAMPLES
The design procedure, presented in Chapter 4, is illustrated
by numerical examples for connections of nonprestressed
slabs with interior, edge, and corner columns. Section D.4 is
a design example of shear reinforcement for a connection of
an interior column with a prestressed slab.
D.1—Interior column-slab connection
The design of headed studs, conforming to ASTM A1044/
A1044M, is required at an interior column (Fig. D.1) based
on the following data: column size cx by cy = 12 × 20 in.2
(305 × 508 mm2
); slab thickness h = 7 in. (178 mm);
concrete cover = 0.75 in. (19 mm); fc′ = 4000 psi (27.6 MPa);
yield strength of studs fyt = 51 ksi (350 MPa); and flexural
reinforcement nominal diameter = 5/8 in. (16 mm). The
factored forces transferred from the column to the slab are:
Vu = 110 kips (489 kN) and Muy = 600 kip-in. (67.8 kN-m).
The five steps of design outlined in Section 4.4 are followed:
Step 1—The effective depth of slab
d = 7 – 0.75 – (5/8) = 5.62 in. (143 mm)
Properties of a critical section at d/2 from column face
shown in Fig. 4.1(a): bo = 86.5 in. (2197 mm); Ac = 486 in.2
(314 × 103
mm2
); Jy = 28.0 × 103
in.4
(11.7 × 109
mm4
); lx1
= 17.62 in. (448 mm); and ly1 = 25.62 in. (651 mm).
The fraction of moment transferred by shear (Eq. (4-3))
γvy 1
1
1
2
3
---
17.62
25.62
-------------+
------------------------------– 0.36= =
Table C.4—Slabs with stud shear reinforcement having s approximately equal to or greater than 0.75d
Exper-
iment
Column size,†
in.
(mm)
fc′,‡
psi (MPa)
d, in.
(mm) s/d
Tested capacities M at critical
section
centroid, kip-
in. (kN-m)
Maximum
shear stress
vu,§
psi (MPa)
fyt, ksi
(MPa)
Av, in.2
(mm2
)
(vu)outside,||
psi (MPa)
vtest/
vcode
**
V,
kips (kN)
M, kip-in.
(kN-m)
(1) (2) (3) (4) (5) (6) (7) (8) (9) (10) (11) (12) (13)
3 11.81 sq. (300 sq.) 3250 (22.4) 8.66 (220) 0.55and
0.73 346 (1540) 0 0 488 (3.37) 47.9 (330) ? 214 (1.48) 1.77
12 7.48 cr. (190 cr.) 4610 (31.8) 6.02 (153) 0.75 159 (707) 0 0 623 (4.30) 67.6 (466) 1.09 (703) 195 (1.34) 1.42
13 7.48 cr. (190 cr.) 3050 (21.0) 5.91 (150) 0.77 128 (569) 0 0 517 (3.66) 57.6 (397) 1.09 (703) 160 (1.10) 1.43
16 9.25 cr. (235 cr.) 4630 (31.9) 5.94 (151) 0.66 207 (921) 0 0 728 (5.02) 72.5 (500) 1.46 (942) 182 (1.26) 1.34
17 9.25 cr. (235 cr.) 5250 (36.2) 6.14 (156) 0.65 216 (961) 0 0 725 (5.00) 72.5 (500) 1.46 (942) 180 (1.24) 1.26
18 9.25 cr. (235 cr.) 5290 (36.5) 6.50 (165) 0.61 234 (1040) 0 0 725 (5.00) 42.5 (293) 1.46 (942) 181 (1.25) 1.26
19 7.87 sq. (200 sq.) 5060 (34.9) 6.30 (160) 0.75 236 (1050) 0 0 661 (4.56) 54.1 (373) 1.40 (903) 165 (1.14) 1.08
21 9.84 sq. (250 sq.) 4100 (28.3) 4.49 (114) 0.70 47.4 (211) 651 (73.6) 491 (55.5) 528 (3.64) 55.1 (380) 0.66 (426) — 1.07
22 9.84 sq. (250 sq.) 4030 (27.8) 4.49 (114) 0.70 52.8 (235) 730 (82.5) 552 (62.4) 590 (4.07) 55.1 (380) 0.66 (426) — 1.20
23 9.84 sq. (250 sq.) 4080 (28.1) 4.49 (114) 0.70 26.9 (120) 798 (90.2) 708 (80.0) 641 (4.42) 55.1 (380) 0.66 (426) — 1.30
24 9.84 sq. (250 sq.) 4470 (30.8) 4.49 (114) 0.70 27.2 (121) 847 (95.7) 755 (85.3) 693 (4.78) 55.1 (380) 0.66 (426) — 1.38
26 9.84 sq. (250 sq.) 4890 (33.7) 4.49 (114) 0.75 33.7 (150) 1434 (162.0) 1434 (162.0) 570 (3.93) 66.7 (460) 1.57 (1010) — 1.02
27 9.84 sq. (250 sq.) 5660 (39.0) 4.49 (114) 0.75 67.4 (300) 1257 (142.0) 1257 (142.0) 641 (4.42) 66.7 (460) 1.57 (1010) — 1.06
30* 9.84 sq. (250 sq.) 5470 (37.7) 4.49 (114) 0.75 117 (520) 0 0 454 (3.13) 40.3 (278) 1.32 (852) — 1.02
31 9.84 sq. (250 sq.) 3340 (23.0) 4.49 (114) 0.75 123 (547) 0 0 476 (3.28) 40.3 (278) 1.32 (852) 136 (0.94) 1.18
32 9.84 sq. (250 sq.) 5950 (41.0) 4.49 (114) 0.75 131 (583) 0 0 509 (3.51) 70.9 (489) 1.32 (852) 145 (1.00) 0.94
33 9.84 sq. (250 sq.) 5800 (40.0) 4.49 (114) 0.75 131 (583) 0 0 509 (3.51) 40.3 (278) 1.32 (852) 145 (1.00) 0.95
34 9.84 sq. (250 sq.) 4210 (29.0) 4.49 (114) 0.75 122 (543) 0 0 473 (3.26) 70.9 (489) 1.32 (852) 166 (1.14) 1.28
35 9.84 sq. (250 sq.) 5080 (35.0) 4.49 (114)
0.75and
1.50
129 (574) 0 0 500 (3.45) 40.3 (278) 1.32 (852) 143 (0.99) 1.00
36 9.84 sq. (250 sq.) 4350 (30.0) 4.49 (114) 0.75 114 (507) 0 0 444 (3.06) 70.9 (489) 1.32 (852) 178 (1.23) 1.35
38 9.84 sq. (250 sq.) 4790 (33.0) 4.49 (114) 0.70 48 (214) 637 (72.0) 476 (53.8) 522 (3.60) 55.1 (380) 0.66 (426) — 1.03
39 9.84 sq. (250 sq.) 4210 (29.0) 4.45 (113) 0.88 113 (503) 0 0 444 (3.06) 47.1 (325) 0.46 (297) — 1.52
40 9.84 sq. (250 sq.) 4240 (29.2) 4.45 (113) 1.00 125 (556) 0 0 492 (3.39) 52.3 (361) 1.74 (1120) 253 (1.74) 1.94
41 9.84 sq. (250 sq.) 5300 (36.6) 4.45 (113) 0.88 133 (592) 0 0 523 (3.61) 49.2 (339) 0.99 (639) 221 (1.52) 1.52
42 9.84 sq. (250 sq.) 5380 (37.1) 4.45 (113) 0.88 133 (592) 0 0 523 (3.61) 49.2 (339) 1.48 (955) 273 (1.88) 1.86
43 12.0 sq. (305 sq.) 4880 (33.7) 4.76 (121) 1.00 134 (596) 0 0 419 (2.89) 73.0 (503) 1.54 (994) 270 (1.86) 1.93
Mean 1.31
*
Slab 30 is semi-lightweight concrete. replaced in calculations by fct/6.7; fct is average splitting tensile strength of lightweight-aggregate concrete; fct used herein = 377 psi (2.60 MPa),
determined experimentally.
†Column 2 gives side dimension of square (sq.) columns, or diameter of circular (cr.) columns.
‡For cube strengths, concrete cylinder strength in Column 3 calculated using fc′ = 0.83fcube′ .
§Column 9 is maximum shear stress at failure in critical section at d/2 from column face.
||(vu)outside in Column 12 is maximum shear stress at failure in critical section at d/2 outside outermost studs; (vu)outside not given for slabs that failed within stud zone.
**vcode is value allowed by ACI 318 in Section 4.3.3. vcode calculated at d/2 from column face when failure is within stud zone and at section at d/2 from outermost studs when failure
is outside shear-reinforced zone.
fc′
--`,``,`,`,`,`,```,`,,``,`,,```-`-`,,`,,`,`,,`---

The maximum shear stress occurs at x = 17.62/2 = 8.81 in.
(224 mm), and its value is (Eq. (4-2))
= 294 psi (2.03 MPa)
(2.70 MPa = 0.52 )
The nominal shear stress that can be resisted without shear
reinforcement at the critical section considered (Eq. (4-7) to
(4-9))
(or 0.37 )
(or 0.38 )
vn = 4 (or /3)
Use the smallest value: vn = 4 = 253 psi ( /3 =
1.74 MPa).
vu
110 10
3
×
486
-----------------------
0.36 600 10
3
×( )8.81
28.0 10
3
×
--------------------------------------------------+=
vu
φ
----
294
0.75
---------- 392 psi = 6.2 fc′= = fc′
vn 2
4
1.67
----------+
⎝ ⎠
⎛ ⎞ fc′ 4.4 fc′= = fc′
vn
40 5.62( )
86.5
--------------------- 2+
⎝ ⎠
⎛ ⎞ fc′ 4.6 fc′= = fc′
fc′ fc′
fc′ fc′
Step 2—The quantity (vu/φ) is greater than vn, indicating
that shear reinforcement is required; the same quantity is less
than the upper limit vn = 8 (in.-lb units) (2 /3 [SI
units]), which means that the slab thickness is adequate.
Stirrups are not permitted by ACI 318 because (vu/φ) is
greater than 6 (in.-lb units) ( /2 [SI units]).
The shear stress resisted by concrete in the presence of
headed studs at the critical section at d/2 from column face
vc = 3 = 190 psi (1.31 MPa)
Use of Eq. (4-1), (4-11), and (4-13) gives
vs ≥ – vc = 392 – 190 = 202 psi (1.39 MPa)
= 0.34 in. (8.7 mm)
Step 3
so ≤ 0.5d = 2.8 in. (71 mm); s ≤ 0.5d = 2.8 in. (71 mm)
This example has been provided for one specific type of
headed shear stud reinforcement, but the approach can be
adapted and used also for other types mentioned in Appendix A.
Try 3/8 in. (9.5 mm) diameter studs welded to a bottom
anchor strip 3/16 x 1 in.2
(5 x 25 mm2
). Taking cover of 3/4 in.
(19 mm) at top and bottom, the specified overall height of
headed stud assembly (having most efficiency) (Eq. (6-1))
ls = 7 – 2 = 5.5 in. (140 mm)
The actual overall height (considering manufacturing
tolerance) should not be less than
ls – 1/2 the diameter of flexural reinforcement bars (5/8 in.)
= 5-3/16 in. (132 mm)
With 10 studs per peripheral line, choose the spacing
between peripheral lines, s = 2.75 in. (70 mm), and the
spacing between column face and first peripheral line, so =
2.25 in. (57 mm) (Fig. D.1)
= 0.40 in. (10.1 mm)
This value is greater than 0.34 in. (8.7 mm), indicating that
the choice of studs and their spacing are adequate.
Step 4—For a first trial, assume a critical section at 4.5d
from column face (Fig. 4.1(b)):
α = 4.5; αd = 4.5(5.62) = 25.3 in. (643 mm);
lx2 = 62.6 in. (1590 mm); ly2 = 70.6 in. (1793 mm); γvy = 0.39
(Eq. (B-2));
fc′ fc′
fc′ fc′
fc′
vu
φ
----
Av
s
-----
vsbo
fyt
----------≥
202 86.5( )
51,000
------------------------=
3
4
---
⎝ ⎠
⎛ ⎞
Av
s
-----
10 0.11( )
2.75
---------------------=
Fig. D.1—Example of interior column-slab connection: stud
arrangement. (Note: 1 in. = 25.4 mm; 1 kip = 4.448 kN.)
--`,``,`,`,`,`,```,`,,``,`,,```-`-`,,`,,`,`,,`---

bo = 209.8 in. (5329 mm);
Ac = 1179 in.2
(760.6 × 103
mm2
); Jy = 547.3 × 103
in.4
(227.8 × 109 mm4).
The maximum shear stress in the critical section occurs on
line AB (Fig. 4.1(a)) at: x = 62.6/2 = 31.3 in. (795 mm);
Eq. (4-2) gives
= 142 psi (0.98 MPa)
The value (vu/φ) = 142 psi (0.98 MPa) is greater than vn =
126 psi (0.87 MPa), which indicates that shear stress should
be checked at α > 4.5. Try 10 peripheral lines of studs; the
distance between column face and outermost peripheral line
of studs is
so + 9s = 2.25 + 9(2.75) = 27 in. (686 mm)
Check shear stress at a critical section at a distance from
column face
αd = 27 + d/2 = 27 + 5.62/2 = 29.8 in. (757 mm)
vu/φ = 125 psi (0.86 MPa)
vn = 2 = 126 psi (0.87 MPa)
Step 5—The value of (vu/φ) is less than vn, which indicates
that the extent of the shear-reinforced zone, shown in Fig. D.1,
is adequate.
The value of Vu used to calculate the maximum shear
stress could have been reduced by the counteracting factored
load on the slab area enclosed by the critical section; this
reduction is ignored in Sections D.2 to D.4.
D.2—Edge column-slab connection
Design the studs required at the edge column-slab connection
in Fig. D.2(a), based on the following data: column cross
section, cx × cy = 18 × 18 in.2
(457 × 457 mm2
); the values
of h, ct, d, fc′, fyt, D, and db, in Section D.1 apply herein. The
connection is designed for gravity loads combined with wind
load in positive or negative x-direction. Cases I and II are
considered, which produce extreme stresses at Points B and
A of the shear-critical section at d/2 from the column or at D
and C of the shear-critical section at d/2 from the outermost
peripheral line of studs (Fig. D.2(a) and (b)). The factored
forces, due to gravity load combined with wind load, are given.
Case I—Wind load in negative x-direction
Vu = 36 kips (160 kN); MuOy = 1720 kip-in. (194 kN-m);
vu
110 10
3
×
1179
-----------------------
0.39 600 10
3
×( )31.3
547.3 10
3
×
--------------------------------------------------+ 107 psi (0.74 MPa)= =
vu
φ
----
107
0.75
----------=
α
29.8
d
----------
29.8
5.62
---------- 5.3= = =
fc′
MuOx = 0
For the shear-critical section at d/2 from column face,
xO = –5.17 in., and Eq. (4-5) gives
Muy = 1720 + 36(–5.17) = 1530 kip-in. (173 kN-m); Mux = 0
Case II—Wind load in positive x-direction
Vu = 10 kips (44 kN); MuOy = –900 kip-in. (–102 kN-m)
Muy = –900 + 10(–5.17) = –952 kip-in. (–107 kN-m)
The five steps of design outlined in Section 4.4 are
followed.
Step 1—Properties of the shear-critical section at d/2 from
column face shown in Fig. D.2(a) are: bo = 65.25 in. (1581 mm);
Ac = 367 in.2
(237 × 103
mm2
); Jy = 17.63 × 103
in.4
(7.338
× 106
mm4
); lx1 = 20.81 in. (529 mm); and ly1 = 23.62 in.
(600 mm).
The fraction of moment transferred by shear (Eq. (B-4))
The shear stress at Points A and B, calculated by Eq. (4-2)
with xA = –14.17 in. or xB = 6.64 in., are given in Table D.1.
The maximum shear stress, in absolute value, occurs at
Point A (Case I) and |(vu/φ)A| = 338/0.75 = 451 psi = 7.1
(3.13 MPa = 0.59 ).
The nominal shear stress that can be resisted without shear
reinforcement at the shear-critical section, vn = 4 = 253 psi
( /3 = 1.74 MPa).
γvy 1
1
1
2
3
---
20.81
23.62
------------- 0.2–+
--------------------------------------------– 0.36= =
fc′
fc′
fc′
fc′
Fig. D.2—Example of edge column-slab connection:
shear-critical sections and stud arrangement. (Note: 1 in.
= 25.4 mm.)
--`,``,`,`,`,`,```,`,,``,`,,```-`-`,,`,,`,`,,`---

Step 2—Because the value (vu/φ) exceeds vn, shear reinforce-
ment is required; the same quantity is less than the upper
limit vn = 8 , psi (2 /3, MPa), indicating that the slab
thickness is adequate.
The shear stress resisted by concrete in presence of headed
studs at the shear-critical section at d/2 from the column face is
vc = 3 = 190 psi ( /4 = 131 MPa)
vs ≥ – vc = 451 – 190 = 261 psi (1.80 MPa)
= 0.33 in. (8.5 mm)
Step 3
so ≤ 0.5d = 2.8 in. (71 mm); s ≤ 0.5d = 2.8 in. (71 mm)
Using 3/8 in. (9.5 mm) diameter studs, arranged as shown
in Fig. D.2(b), with so = 2.25 in. (57 mm) and s = 2.75 in.
(70 mm) gives: (Av/s) = 9(0.11)/2.75 = 0.36 in. (9.1 mm).
This value is greater than 0.33 in. (8.5 mm), indicating that
Step 4—Try nine peripheral lines of studs; the properties
of the shear-critical section at d/2 from the outermost peripheral
line of studs are:
bo = 132 in. (3353 mm); Ac = 742 in.2 (479 × 103 mm2); Jy
= 142.9 × 103 in.4 (59.48 × 109 mm4);
lx2 = 45 in. (1143 mm); ly2 = 72 in. (1829 mm); γvy = 0.30
(Eq. (B-4));
xC = –27.6 in. (–701 mm); xD = 17.4 in. (445 mm); xO =
–18.6 in. (–472 mm).
The factored shearing force and unbalanced moment at an
axis, passing through the centroid of the shear-critical
section outside the shear-reinforced zone, are (Eq. (4-5)):
Case I: Vu = 36 kips (160 kN); Muy = 1720 + 36(–18.6)
= 1050 kip-in. (118 kN-m)
Case II: Vu = 10 kips (44 kN); Muy = –900 + 10(–18.6)
= –1090 kip-in. (–123 kN-m)
fc′ fc′
fc′ fc′
vu
φ
----
Av
s
-----
vsbo
fyt
----------≥
261 65.25( )
51,000
---------------------------=
Equation (4-2) gives the shear stresses at Points C and D,
listed in Table D.1 for Cases I and II.
Point D (Case I) and |(vu/φ)D| = 87/0.75 = 116 psi = 1.8
(0.80 MPa = 0.15 ). The nominal shear strength outside
the shear-reinforced zone, vn = 2 = 126 psi (0.17 =
0.87 MPa).
Step 5—The value of (vu/φ) is less than vn, indicating that
the extent of the shear-reinforced zone, as shown in Fig. D.2(b),
is adequate.
D.3—Corner column-slab connection
The corner column-slab connection in Fig. D.3(a) is
designed for gravity loads combined with wind load in positive
or negative x-direction. The cross-sectional dimensions of
the column are cx = cy = 20 in. (508 mm) (Fig. D.3(a)). The
same values of: h, ct, d, fc′ , fyt, D, and db, in Section D.1
apply in this example. Two cases (I and II) are considered,
producing extreme shear stresses at Points A and B of the
shear-critical section at d/2 from the column or at C and D of
the shear-critical section at d/2 from the outermost peripheral
line of studs (Fig. D.3(a) and (b)). The factored forces, due
to gravity loads combined with wind load, are given.
Case I—Wind load in positive x-direction
Vu = 6 kips (27 kN); MuOy = –338 kip-in. (–38 kN-m);
MuOx = 238 kip-in. (27 kN-m)
For the shear-critical section at d/2 from column face, xO
= yO = –7.11 in. (–181 mm) and θ = 45 degrees; thus, Eq. (4-5)
and (4-6) give
Muy = –338 + 6(–7.11) = –381 kip-in.;
Mux = 238 + 6(–7.11) = 195 kip-in.
Muy = –132 kip-in. (–15 kN-m);
Mux = 407 kip-in. (46 kN-m)
fc′
fc′
fc′ fc′
Table D.1—Shear stresses* (psi) due to factored
loads; edge column-slab connection (Fig. D.2)
Shear-critical section Case I Case II
At d/2 from column face
(vu)A (vu)B (vu)A (vu)B
–338 302 299 –100
At d/2 from outermost peripheral line
of studs
(vu)C (vu)D (vu)C (vu)D
–13 87 77 –27
*
vu represents stress exerted by column on slab, with positive sign indicating upward
stress.
Note: 1 MPa = 145 psi.
Fig. D.3—Example of corner column-slab connection:
shear-critical sections and stud arrangement. (Note: 1 in. =
25.4 mm.)
--`,``,`,`,`,`,```,`,,``,`,,```-`-`,,`,,`,`,,`---

Case II—Wind load in negative x-direction
Vu = 22 kips (97 kN); MuOy = 953 kip-in. (108 kN-m);
MuOx = 377 kip-in. (43 kN-m)
Muy = 953 + 22(–7.11) = 797 kip-in.;
Mux = 377 + 22(–7.11) = 221 kip-in.
Muy = 720 kip-in. (81 kN-m);
Mux = –407 kip-in. (–46 kN-m)
The five steps of design, outlined in Section 4.4, are followed.
Step 1—Properties of the shear-critical section in Fig. D.3(a)
are: bo = 45.63 in. (1159 mm); Ac = 257 in.2
(166 × 103
mm2
);
Jx = 22.26 × 103 in.4 (9.27 × 109 mm4) and Jy = 5.57 × 103 in.4
(2.32 × 109
mm4
). The projections of the critical section on
the x and y axes are: lx1 = 16.13 in. (410 mm); and ly1 =
32.26 in. (820 mm). The fractions of unbalanced moments
transferred by shear are (Eq. (B-5) and (B-6))
; γvx = 0.4
The factored shear stress at Point A (–8.07, 16.13 in.) in
Case I is (Eq. (4-2))
= 192 psi (1.33 MPa)
Similar calculations give the values of vu at Points A and B
(8.07, 0 in.) for Cases I and II, which are listed in Table D.2.
Point B (Case II) and |(vu/φ)B| = 364/0.75 = 485 psi =
7.7 (3.35 MPa = 0.64 ). The nominal shear stress
that can be resisted without shear reinforcement at the shear-
critical section, vn = 4 = 253 psi ( /3 = 1.74 MPa)
(Eq. (4-7) to (4-9)).
Step 2—Because the value (vu/φ) exceeds vn, shear
reinforcement is required; the same quantity is less than the
upper limit, vn = 8 (in.-lb units) (2 /3 [SI units]),
indicating that the slab thickness is adequate.
headed studs at the shear-critical section at d/2 from the
column face is
vc = 3 = 190 psi ( /4 = 1.31 MPa)
Use of Eq. (4-1), (4-11) and (4-13) gives
vs ≥ – vc = 485 – 190 = 295 psi (2.03 MPa)
= 0.26 in. (6.7 mm)
γvy 1
1
1 2/3( ) lx1/ly1( ) 0.2–+
-----------------------------------------------------------– 0.267= =
vu( )A
6 10
3
×
257
-----------------
0.4 407 10
3
×( )16.13
22.26 10
3
×
--------------------------------------------------
0.267 132 10
3
×–( ) 8.07–( )
5.57 10
3
×
----------------------------------------------------------------+ +=
fc′ fc′
fc′ fc′
fc′ fc′
fc′ fc′
vu
φ
----
Av
s
-----
vsbo
fyt
----------≥
295 45.63( )
51,000
---------------------------=
Step 3
so ≤ 0.5d = 2.8 in. (71 mm); s ≤ 0.5d = 2.8 in. (71 mm)
Using 3/8 in. (9.5 mm) diameter studs, arranged as shown
in Fig. D.3(b), with so = 2.25 in. (57 mm) and s = 2.5 in.
(64 mm) gives: (Av/s) = 6(0.11)/2.5 = 0.26 in. (6.7 mm). This
value is the same as that calculated in Step 2, indicating that
Step 4—Try seven peripheral lines of studs; the properties
of the shear-critical section at d/2 from the outermost peripheral
line of studs (Fig. D.3(b)) are:
xO = yO = –17.37 in. (–441 mm); θ = 45 degrees;
bo = 69 in. (1754 mm); Ac = 388 in.2
(251 × 103
mm2
);
Jx = 116.9 × 103
in.4
(48.64 × 109
mm4
); Jy = 9.60 × 103
in.4
(4.00 × 109
mm4
);
lx2 = 15.0 in. (380 mm); ly2 = 56.7 in. (1439 mm); γvx = 0.40
(Eq. (B-5)); γvy = 0.14 (Eq. (B-6)).
The factored shearing force and unbalanced moment about
the centroidal principal axes of the shear-critical section
outside the shear-reinforced zone (Eq. (4-5) and (4-6)), are:
Case I:
Vu = 6 kips (27 kN); Mux = 407 kip-in. (46 kN-m);
Muy = –218 kip-in. (–25 kN-m)
Case II:
Vu = 22 kips (97 kN); Mux = –407 kip-in. (–46 kN-m);
Muy = 402 kip-in. (45 kN-m)
Use of Eq. (4-2) gives the values of vu at Points C (–10.38,
28.33 in.) and D (4.59, 13.36 in.) for Cases I and II, listed in
Table D.2.
Point C (Case I) and |(vu/φ)C| = 89/0.75 = 119 psi = 1.88
(0.82 MPa = 0.16 ). The nominal shear stress outside the
shear-reinforced zone, vn = 2 = 126 psi (0.17 =
0.87 MPa).
Step 5—The value of (vu/φ) is less than vn, indicating that
the extent of the shear-reinforced zone, as shown in Fig. D.3(b),
is sufficient.
D.4—Prestressed slab-column connection
Design the shear reinforcement required for an interior
column, transferring Vu = 110 kips (490 kN) combined with
fc′
fc′
fc′ fc′
Table D.2—Shear stresses* (psi) due to factored
loads; corner column-slab connection (Fig. D.3)
Shear-critical section Case I Case II
At d/2 from column face
(vu)A (vu)B (vu)A (vu)B
192 –28 –312 364
At d/2 from outermost peripheral line
of studs
(vu)C (vu)D (vu)C (vu)D
89 19 –46 65
*
vu represents stress exerted by column on slab, with positive sign indicating upward
stress.
Note: 1 MPa = 145 psi.
--`,``,`,`,`,`,```,`,,``,`,,```-`-`,,`,,`,`,,`---

unbalanced moment Muy = 550 kip-in. (62 kN-m) to a post-
tensioned flat plate of thickness, h = 7 in. (178 mm). The slab
has equal spans 280 x 280 in.2 (7.1 x 7.1 m2). The column
size is 16 x 16 in.2
(406 x 406 mm2
). The values of ct, fc′, fyt,
and D, in Section D.1 apply herein. Tendon profiles are
commonly composed of parabolic segments, for which the
average effective prestress fpc, required to balance a fraction
κ of the self-weight, (hγconc) per unit area, plus the superim-
posed dead load of intensity wsd can be calculated as (Gayed
and Ghali 2006) (Fig. D.4(a))
(D-1)
where L is the panel length, and geometrical parameters: α
and hc are defined in Fig. D.4(a). Choose the values: κ =
0.85; α = 0.1; γconc = 153 lb/ft3 (24 kN/m3); wsd = 27 lb/ft2
(1.3 kPa); L = 280 in. (7.1 m); h = 7 in.; and hc = 3.8 in.
Equation (D-1) gives fpc = 202 psi (1.39 MPa). This level of
prestressing is closely acquired by ten 0.6 in. seven-wire
post-tensioned nonbonded strands per panel. The cross-
sectional area per strand = 0.217 in.2
(140 mm2
); the average
value of the effective compressive stress provided by ten
tendons in each of two directions is
= 194 psi (1.34 MPa)
fpc
κ 1 2α–( ) γconch wsd+( )L
2
8hhc
-----------------------------------------------------------------=
fpc
10 38 10
3
×( )
280 7( )
-------------------------------=
Prestressing tendons are typically placed in bands over
support lines in one direction and uniformly distributed in
the perpendicular direction. In the current example, the
prestressing tendons are banded in the x-direction and
uniformly distributed in the y-direction (Fig. D.4(b)). ACI
318 requires that at least two tendons should pass through the
column cage in each direction; the arrangement of the
tendons as shown in Fig. D.4(b) satisfies this requirement.
ACI 318 requires a minimum amount of bonded top flexural
reinforcing bars in the vicinity of the column; choose eight
bars of diameter db = 1/2 in.; for clarity, the bonded bars are
not shown in Fig. D.4. A check that the cross-sectional areas
of the bonded and nonbonded reinforcements satisfy the
ultimate flexural strength required is necessary, but is
beyond the scope of the present report.
Punching shear design: Vu = 110 kips (490 kN); Muy =
550 kip-in. (62 kN-m).
The five steps of design, outlined in Section 4.4, are followed.
Step 1—Properties of the shear-critical section at d/2 from
the column are: d = h – ct – db = 7 – 3/4 – 1/2 = 5.75 in.; bo
= 87 in. (2210 mm); Ac = 500 in.2 (323 × 103 mm2); Jy = 39.4
× 103 in.4 (16.4 × 109 mm4); lx1 = ly1 = 21.75 in. (552 mm);
and γvy = 0.4 (Eq. (B-2)). The maximum shear stress occurs
at x = 21.75/2 = 10.88 in. (276 mm), and its value is (Eq. (4-2))
= 281 psi (1.94 MPa)
= = 375 psi = 5.9 (2.59 MPa = 0.49 )
The three conditions, warranting the use of Eq. (5-1a) or
(5-1b), are satisfied at the considered connection. Two tendons
from each direction intercept the critical section at d/2 from
the column; the sum of the vertical components of these
tendons at the location of the shear-critical section, Vp = 6 kips
(26 kN). It is uncertain that the actual cable profiles, in the x
and y directions, will have slopes matching those used in
calculating Vp (≈ 0.02); thus, for safety, assume that Vp = 0.
Substituting the values of fpc and Vp in Eq. (5-1a) gives
vn = 3.5 + 0.3(194) + 0 = 280 psi (1.93 MPa)
Step 2—The quantity (vu/φ) is greater than vn, indicating
that shear reinforcement is required; the same quantity is less
than the upper limit vn = 8 (in-lb units) (2 /3 [SI
units]), which means that the slab thickness is adequate.
headed studs at the critical section at d/2 from column face
vc = 3 = 190 psi ( /4 = 1.31 MPa)
vs ≥ – vc = 375 – 190 = 185 psi (1.28 MPa)
vu
110 10
3
×
500
-----------------------
0.40 500 10
3
×( )10.88
39.4 10
3
×
-----------------------------------------------------+=
vu
φ
----
281
0.75
---------- fc′ fc′
4000
fc′ fc′
fc′ fc′
vu
φ
----
Fig. D.4—Example connection of interior column-prestressed
slab. (Note: 1 in. = 25.4 mm.)
--`,``,`,`,`,`,```,`,,``,`,,```-`-`,,`,,`,`,,`---

= 0.32 in. (8.0 mm)
Step 3—(vu/φ) < 6 (psi); thus, stirrups or headed studs
can be used. For ease of installation of the prestressing
tendons, use studs with s ≤ 0.75d. Because the column width
is large with respect to d, eight studs per peripheral line will
not satisfy the requirement g ≤ 2d (Fig. 1.2); choose 12 studs
per peripheral line.
so ≤ 0.5d = 2-7/8 in. (73 mm); s ≤ 0.75d = 4-5/8 in. (117 mm)
With twelve 3/8 in. studs per peripheral line and spacing s
= 4 in. (102 mm),
= 0.33 in. (8.4 mm)
This value is greater than 0.32 in., indicating that the
choice of studs and their spacing are adequate.
Step 4—Try seven peripheral lines of studs. Properties of
critical section at d/2 from the outermost peripheral line of
studs (Fig. D.4(b)) are:
lx2 = ly2 = 75.5 in.; γvy = 0.4 (Eq. (B-2)); bo = 235 in.; Ac
= 1351 in.2
; and Jy = 848.2 × 103
in.4
.
The maximum shear stress in the critical section occurs at:
x = 75.5/2 = 37.8 in. (959 mm); Eq. (4-2) gives
= 91 psi (0.63 MPa)
vn = 2 = 126 psi (0.17 = 0.87 MPa)
Step 5—The value of (vu/φ) = 91/0.75 = 121 < 126 psi,
indicating that the extent of the shear-reinforced zone, as
shown in Fig. D.4(b), is adequate.
Av
s
-----
vsbo
fyt
----------≥
185 87( )
51,000
--------------------=
fc′
Av
s
-----
12 0.11( )
4
---------------------=
vu
110 10
3
×
1351
-----------------------
0.40 550 10
3
×( )37.8
848.2 10
3
×
--------------------------------------------------+=
fc′ fc′
--`,``,`,`,`,`,```,`,,``,`,,```-`-`,,`,,`,`,,`---

Aci 421-1 r-08

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Aci 421-1 r-08